Physicochemical Aspects of

Steelmaking Processes*

Reactions in Ironmaking and

By E. T. T URKDOGAN**

Synopsis This is a critical review of the physical chemistry of selected aspects of ironmaking and steelmaking processes. The equilibrium data from gas-slag-metal reactions are presented in a form suitable for easy application in the study of the state of reactions in ironmaking and steelmaking. Plant data indicate departures from equilibrium for most reactions in the blast-

furnace bosh and hearth. In steelmaking, most reactions are close to equi-librium at the time of furnace tapping. Comments are made on reaction equilibria in relation to the ladle refining of hot metal. A summary is made of new equilibrium data for desulfurization of liquid steel with cal-cium aluminate slags. Various examples are given of the role of vapor species in pyrometallurgical processes, e.g., silicon- and sulfur-bearing spe-cies and alkali cyanide vapors in the blast furnace, causes of fume emission during furnace tapping, and consequences of fume emission in coal gasifica-tion using liquid iron as the reaction medium.

I. Introduction

Studies of the physical chemistry of gas-slag-metal reactions and measurements of the physicochemical and thermochemical properties of materials pertinent to ironmaking and steelmaking processes have gone through many interesting stages of evolution during the past six decades or more. The first textbook writ-ten on the physical chemistry of steelmaking was the classical work by Schenck, " Einfuhrung in die Phys-ikalische Chemie der Eisenhuttenprozesse," published in the years 1932 and 1934. Belated English transla-tion of this book in 19451) was of much value to steel-makers in many countries and stimulated the birth of a new era of research on the physical chemistry of metallurgical reactions. The symposium on physical chemistry organized by the Faraday Society in 19482) set the main theme of research for the subsequent two decades on the fundamentals of pyrometallurgical pro-cesses. Accumulated knowledge of the physical chemistry of gas-slag-metal reactions and the physicochemical

properties of solid and liquid metallic solutions, mol-ten slaps, and refractory materials have been critically assessed and compiled by Kubaschewski and Alcock,3~ Richardson,4) and Turkdogan.5's) These reference books are used for the background scientific and tech-nical information in the preparation of the present

paper. Upon kind invitation by the Iron and Steel Institute of Japan, a critical assessment is presented in this pa-

per of some aspects of the present state of knowledge of the physicochemical aspects of ironmaking and steelmaking processes. For the paper to be technical-ly comprehensive, the subject of discussion is confined to a few selected topics of process metallurgy in iron-making and steelmaking that are of current interest:

(1) the state of reactions in the blast furnace and the converter in relation to the gas-slag-metal reaction equilibria; (2) ladle refining; and (3) the role of vapor species in metallurgical reactions, including coal

gasification in liquid iron bath.

II. Gas-Slag-Metal Reaction Equilibria

1. General Considerations Gas-slag and metal-slag reactions are redox reac-

tions which are accompanied by exchange of electrons between the reacting species as they are transferred to and from nonpolar gaseous or metallic phase and ionic molten slag. For example, the transfer of sulfur from metal to slag is a reduction reaction involving ex-change of electrons between sulfur and oxygen atoms:

5-E-(O2-) = 0+(S2) ..................(1)

where underscore indicates solute in solution in the metal and ( ) species dissolved in the slag. In most cases, the transfer of a reactant from metal to slag is an oxidation process, such as

P+ 5 0 + 3 (02-) = (POr) ...............(2) 2 2

Although the reactions as formulated above are conceptually correct, their equilibrium constants can-not be evaluated because the thermodynamic activity or activity coefficient of an individual ion cannot be determined experimentally. However, numerous attempts have been made in the past, following the work of Herasymenko,7'8) to represent the equilib-rium constants of slag-metal reactions in terms of assumed ionic concentrations.9'1°) Although the deri-vation of reaction equilibrium relations from the ionic structural models are of theoretical interest, their ap-

plication to gas-slag-metal reaction equilibria in mul= ticomponent systems has not been particularly reward-ing when compared with the conventional methods of representing the reaction equilibria in terms' of the measured thermodynamic quantities.

The solution models of slaps, either constitutional or structural, have been much exploited in pursuit of a rationale for the composition dependence of the ac-tivity of constituent oxides in slaps. No matter what major differences there might be between assumed solution models, they all yield activity vs, composition relations in reasonable agreement with the same set of data for binary and ternary silicate melts, because of the empirically adjustable parameters involved in

*

**

Received November 14, 1983.

Technical Center, United States Steel Corporation, Monroeville, PA 15146, U.S.A.

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(592) Transactions ISIJ, Vol. 24, 1984

the formulations. However, although the solution

models may be fictitious, the formalisms derived there-from are helpful in the extrapolation of data to other composition ranges. For example, Ban-ya et a1.11'12)

have shown that the composition dependence of the

activity coefficient of ferrous oxide in complex silicate and phosphate melts can be represented by an em-

pirical formalism based on the regular solution model for polymeric melts proposed by Lumsden.l3~

Because of lack of activity data for complex slaps and limited applicability of empirical solution models,

the slag-metal equilibrium relations are often repre-sented in terms of the concentrations of reactants and

products by adopting the mass-action law. The iso-thermal equilibrium constant K thus expressed varies

with the slag composition, on account of omission of the activity coefficients of oxide reactants and prod-

ucts in the expression of the equilibrium constant K. The mass-action law has been applied in the past

in a variety of ways to represent the state of slag-

metal reaction equilibria. The apparent disparities sometimes seen in the experimental data of different

investigators on slag-metal equilibria are often at-tributable to the format of representing the state of

equilibrium for slags of varying compositions. In an attempt to develop a relatively simple rationale, Turk-

dogans~ proposed a method of application of phase rule to describe the slag-metal reaction equilibria as

discussed below. Let us consider in a general way the equilibrium

distribution of an element X between slag and metal.

X (in metal) _ (X) (in slag) (3)

The independent state variables which determine the solubility of X in the slag (in a simple or complex ionic form) coexistent with the metallic phase may be deduced by invoking the Gibbs phase rule:

p+f =c+2 (4)

where, p : the number of phases f : the number of degrees of freedom

c: the number of components of the system. For the phosphorus reaction between liquid iron and slag, the system may be described by the following six components: Fe, P, 0, (CaO), (SiO2) and (MO), where MO represents MgO plus other minor slag components, e.g., MnO, A12O3, CaF2, ... For isobaric and isothermal conditions, we now have from Eq. (4)

p+ f = 6. With the assumption that phosphorus has virtually no effect on its activity coefficient in the metal and slag phase, the equilibrium distribution ratio (%P)/[%P] between slag and metal may be de-scribed by the following functional relations, depend-ing on the number of phases present* : 2 phases : slag and metal,

L [%P] 1p,T(5)

where, V

3 phases:

: the basicity.

MgO-saturated slag and metal,

(%P)

4 phases

[%P] P,T{[%O] , (%SiO2), V} (6)

: MgO- and CaO-saturated slag and metal,

[(%P) [[%P] P,T

~b{[%o] , (%5102)} (7)

Although qualitative, the slag basicity defined by the concentration ratio (oxides of network modifiers)/ (oxides of network formers) gives a general indication of the extent of depolymerization of the melt, hence is a simple indicator of the trend in changes of oxide activities with composition. Consequently, the equi-librium constants for slag-metal reactions expressed in terms of concentrations of slag components are often simple functions of slag basicity, at least for composi-tion ranges of practical interest.

The concept of slag basicity appears to have been first introduced by Herty for steelmaking slags in the mid 1920's; he proposed a simple concentration ratio, so-called the V ratio:

%CaO

The definition of slag basicity is somewhat arbitrary. For example, if we were to assume that in steelmaking-type slags the concentrations of CaO and MgO are equivalent on a molar basis, and similarly on a molar basis 1 /2 P2O5 is equivalent to SiO2j in terms of mass concentrations of oxides, the slag basicity B may be defined by the ratio:

B = %CaO+1.4(%MgO) (9) ...............

%SiO2+0.84(%P2O5)

The basic steelmaking slags are saturated with mag-nesia and dicalcium (magnesium) silicate. For such slaps containing less than about 3 % P2O5, the basicity B given by the above expression is essentially a linear function of the V ratio with a proportionality factor of about 1.17.

B =1.17 %CaO .....................(10) %5102

For blast furnace slags, which contain 10 to 20 % A12O3, the basicity is often represented by the mass concentration ratio:

_ %CaO + %MgO 11

2. Ironmaking Reactions

The state of equilibrium in metallurgical reactions involving three or more phases cannot be evaluated

accurately without giving due consideration to inter-relations among different reactions occurring simul-

taneously. In conventional laboratory experiments

with blast-furnace-type systems at unit activities of carbon and carbon monoxide, many hours of reaction

* It should be noted that the argument presented here is the corrected version of that given by the author in Ref . 6).

Review

time are needed to reach the equilibrium distribution of silicon, manganese, and sulfur between slag and metal to satisfy the three-phase equilibria in the over-all reactions:

(Si02)+2C = Si+2C0 .....................(12)

(CaO)+S+C = (CaS)+CO ............(14)

The studies made by Turkdogan et a1.14) have re-vealed that a partial equilibrium in such three-phase systems could be realized within about 1-hr reaction time using small amounts of slag and metal samples. They obtained the equilibrium relations for the fol-lowing three coupled reactions in graphite-saturated melts and blast-furnace-type slags containing 10 to 20 % A1203.

2(MnO)+Si = (Si02)+2Mn ...............(15)

(CaO)+ 1 Si+S = (CaS)+ 1(SiO2) ......(16) 2 2

(CaO)+Mn+S = (CaS)+(MnO) .........(17)

From their own equilibrium measurements and those of Filer and Darken,15) Turkdogan et a1.14) de-rived the following equilibrium relation for reaction

(15)

[%Mn] 2 (%Si02) KM11s1= (%Mn0) [%Si]

log KMl,Si = 2.8B'-l.16

where, B': 1 400 Within the

effect of to Togethe

Darken,15) equilibrium Stobo,17 ti

was represented by

(%S) (%SiO2) lie 1 Kb S - -[%SJ - [%Si} ) (%C

log K515 = 6327 -4.43+1.4B'

the slag basicity as given by Eq. (11) . to 1 600 °C investigated, the range

mperature on KMnsl could not be detected. with the experimental data of Filer and r

Hatch and Chipman,ls) and Taylor and relation for reaction (16)

e

a0)

(19)14)

Equations (18) and (19) give the following equi-librium relation for the manganese-sulfur coupled re-action (17).

(%S) (%MnO) 1 'Mils = -[%S] [Mn] (%Ca0) 20 6327 log KMfS = T _3

- .85

In most blast furnace operations, the sum of the concentrations of MgO, MnO, and A1203 (=MO) does not change much. Also, in melts saturated with graphite in equilibrium with a fixed partial pressure of carbon monoxide, the activity of oxygen remains constant for a given temperature. For these limiting conditions, the functional relation for the slag-metal equilibrium is simplified to

Transactions ISIJ, Vol. 24, 1984 (593)

(%X) _ [%X] 0 {B'} (21)

The equilibrium distributions of manganese and silicon between slag and metal shown by dotted curves in Fig. 1 are consistent with the functional relation

(21) deduced from the general thermodynamic con-siderations. Since CaO is one of the reactants as

given by Eq. (14), the sulfur distribution ratio has to be a function of both the slag basicity B' and the con-centration of CaO in the slag; the dotted curves in Fig. 1 for (%S)/[%S] are for 40 % CaO in the slag. The dotted equilibrium curves in Fig. 1 for graphite-saturated melts at 1 bar CO pressure were derivedl4~ from the appropriate free energy data, oxide activities, and experimentally determined equilibrium relations

given by Eqs. (18) and (19). In his 1978 Howe Memorial Lecture, Turkdoganls)

discussed the state of slag-metal reactions in the blast furnace hearth, based on the analysis of the daily aver-age plant data for various blast furnaces of U.S. Steel Corporation and those reported by Okabe et al.l 9> for Chiba Works of Kawasaki Steel Corporation. The data analyzed are typical for most blast furnace opera-tions. The melt temperatures at tap are in the range 1 450 to 1 550 °C, mostly in the range 1 475 to 1525 °C. The average slag compositions in mass percent are in the range : 38 to 44 % CaO, 8 to 10 % MgO, 34 to 38 % Si02, 10 to 12 % A1203, 0.5 to 1.0 % MnO, 1 to 2 % S, 0.1 to 0.6 % K20, <0.2 % FeO, and minor amounts of other oxides. The sum of the concentrations of four major constituent oxides CaO, MgO, Si02, and A1203 is in the range 95 to 97 %. The basicity of these slags, defined by the ratio (% CaO+% MgO)/% Si02, is in the range 1.25 to 1.55. The hot metal compositions from these blast furnaces are in the range : 0.4 to 0.8 % Mn, 0.5 to 1.5 % Si, 0.02 to 0.05 % S and other usual impurities, the amount of carbon (%/5 %) in graphite-saturated metal depending on temperature and the concentra-tion of the other solutes.

The plant data are scattered, irrespective of the hot metal temperature at tap, within the shaded areas shown in Fig. 1 for silicon, manganese, and sulfur dis-tribution ratios. However, for a given blast furnace, the scatter in the data is usually much less and the effect of slag basicity on the distribution ratios is quali-tatively in accord with the equilibrium data. That is, for any given blast furnace, the daily average plant data give the ratio [%Si]/(%S102) that decreases with increasing slag basicity; [%Mn]/(%MnO) and (%S)/

[%S] increase with increasing slag basicity. Of the three reactions considered, the distribution of silicon between metal and slag is closer to equilibrium for the three-phase gas-slag-metal reaction (12) in the blast furnace hearth.

If there is approach to equilibrium for the man-ganese-sulfur coupled reaction (17), for small changes in the concentration of CaO in slags, the sulfur-dis-tribution ratio should be directly proportional to the manganese-distribution ratio. The plant data are scat-tered, irrespective of temperature, within the shaded

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(594) Transactions ISIJ, Vol. 24, 1984

area shown in Fig. 2 (top diagram) ; the equilibrium lines (dotted) are from Fq. (20)) for 40 % CaO. How-ever, for a given blast furnace there is much less scatter in the data which show the expected linear relation between (%S)/[%S] and [%Mn]/(%MnO).

Values of KMnsi from plant data scattered within the shaded area in Fig. 2 (bottom diagram) are be-low the equilibrium line (dotted). This observation from the plant data may indicate that as the metal droplets pass through the slag layer in the blast fur-nace, the direction of the reaction might be

2(MnO)+Si -> 2Mn+(Si02) ............(22)

Such an intuitive deduction, however, may not be cor-rect. Observed departure from equilibrium for re-action (15) may be attributed to competition between FeO and MnO in the slag for oxidation of silicon. It is a generally accepted view that with a blast-furnace burden of low basicity, the residual iron oxide in the melting zone of the bosh is higher than with a burden of high basicity. In slags of low basicity, the higher concentration of iron oxide in the upper part of the slag layer would be in competition with MnO for re-action with both silicon and carbon. Consequently, departure from equilibrium for reactions (13) and (14) is expected to be greater in slags of low basicity. The trend seeing from the plant data in Figs. 1 and 2 sub-stantiate the deductions made from the foregoing argu-ment. A study of plant data from a ferromanganese

blast furnace20~ also show departures from equilibrium for reaction (15) in a manner similar to that for the iron blast furnace.

Delve et a1.21~ have studied the titanium-silicon re-action in graphite-saturated melts using blast-furnace-type slags

(TiO2)+Si = (SiO2)+Ti ...............(23)

Their experimental results, for temperatures of 1 500 and 1 600 °C and basicities of 1 to 2, may be sum-marized by the following equation.

KTiSi = [%Ti] _ (%S102) (%Ti02) [%SiJ 24

%Ca0 log KTisi = 0.46 - - - %Si02 )+0.39 From the blast-furnace-plant data compiled by Delve et al.21j and by Hess et al.,22~ it is noted that the values of KTisi are two to three times lower than those given by Fq. (24), which presumably represents the equi-librium relation. Observed departure from equilib-rium for reaction (23) in the blast furnace hearth is similar to the state of silicon-manganese reaction (15) in iron- and ferromanganese-blast furnaces.

3. Steelmaking Reactions

Many equilibrium measurements have been made of reactions between liquid iron alloys and slaps rele-vant to steelmaking processes. The equilibrium rela-

Fig. 1. Daily average blast furnace data are compared with equilibrium relations for slag/metal distribution of silicon, manganese and sulfur in melts saturated with graphite at 1 bar pressure of CO.

Review

Transactions Isu, Vol. 24, 1984 (595)

tions for the key reactions given below are taken from the author's critical assessment6r of the experimental data of various origin.

The recommended value of the equilibrium con-stant for the carbon-oxygen reaction in liquid iron is that obtained by Fuwa and Chipman23} :

C0=C+0

K = [ac] [ao] Pco ............(25)

1 168 log K = - T --2.076

where Pco is in bar and the activities ac and ao are taken equivalent to [%C] and [%O], respectively, at low solute concentrations. It should be noted that the ratio C02/CO in the gas increases with decreasing concentration of carbon. For a total pressure of Pco2+Pco=1 bar, the CO2+C=2C0 equilibrium

gives for 1 600 °C : 0.05 bar CO2 for 0.03 % C and 0.14 bar CO2 for 0.01 % C in iron.

From the known oxygen solubility in liquid iron and the iron oxide activities in the quaternary melts CaO-MgO-FeO~-SiO2, the following relation is ob-tained for CaO (MgO)-saturated simple stags in equi-librium with iron-carbon melts (< 1 % C) at 1 600 °C and CO

2+CO=1 bar.

[%C](%FeOt)-0.33 (26)

where the subscript t indicates the total iron in the slag as oxides represented as FeO, i.e., %FeOt=

%FeO+0.9 X %Fe2O3. For the slag basicity V> 1, the equilibrium relation

for the manganese reaction is reduced to a simple form for temperatures of 1 550 to 1 700 °C.

Mn+(FeO) = (MnO)+Fe

KMn~e = (%MnO) [%Mn](%FeOt)

%CaO - - %SiO2 )KMnFe = 6

F

1 d h saturated with

(27)

or s ags CaO and Ca4P2O9, the

following equilibrium relation was obtained by Ban-

ya and Matoba24~ for the phosphorus oxidation at low concentrations in iron.

4CaO(s)+2P+50 = Ca4P2O9(s) 1 K _ P [%P]2[~/pO]5 ......(28)

log Ky = 75- 5 660660-31.77

The temperature dependence of KP determined in an

earlier study by Fischer and vom Fnde25~ differs con-siderably from the above equation. However, for the

temperature range 1 530 to 1 585 °C investigated by Ban-ya and Matoba,24~ the values of KP in these in-

Fig. 2.

Daily average blast furnace data are compared with

equilibrium relations for manganese-sulfur and

manganese-silicon coupled reactions in melts satu-

rated with graphite.

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[596) Transactions ISIJ, Vol. 24, 1984

dependent determinations agree within a factor of about 1.5. It should be noted that the enthalpy change accompanying reaction (28) estimated from the thermochemical data is closer to that given by the above equation.

Various formalisms have been used in the past to represent the slag-metal equilibrium with respect to the phosphorus reaction."26-29) Using the slag-metal equilibrium data and thermochemical data available in the early 1950's, Pearson and Turkdogan30) derived an empirical equation for the effect of temperature and slag composition on the activity coefficient of P205 in the slag, relative to the hypothetical pure liquid P205. A similar formalism was adopted by Suito et a1.31) to interpret their recent results for phos-phorus distribution between liquid iron and slags satu-rated with magnesia. These different formalisms of varying complexity do not fit all the experimental data on the phosphorus reaction.

A new formalism proposed recently by the author° is based on the functional relations given by Eqs. (5) to (7). Using all the available experimental data, including the recent work of Suito et al.,31) it was found that for slaps not saturated with lime or dical-cium silicate, the following functional relation de-scribed the data reasonably well*.

(%P205) KPS = - -~%P~ - (1-%Si02) = {%FeOt, V}

........................(29)

For double saturation of the slag with lime and mag-nesia, when the basicity V>2.5, the isothermal equi-librium relation KPS is a function % FeOt and essen-tially independent of basicity, in accord with the derivation of the functional relation (7).

For slags of the system CaO-FeO-Si02 P205 satu-rated with either CaO or Ca4P209, the phosphorus distribution ratio would be a function of both % FeO~

and % Si02. However, the data of Fischer and vom Ende25) and of Knuppel and Oeters28) in Fig. 3 for these singly saturated phosphate melts may be repre-sented by a single curve. The scatter in the data masks the apparently small effect of % Si02 on the relation:

(%P205) -'"°'F eO ~} ...............(30) -- [%P}

for singly saturated slags containing less than 8 % S102. From the analysis of available experimental data, the relations in Fig. 4 were obtained for steelmaking-type slags saturated with both dicalcium silicate and magnesio-wustite, for general guidance in evaluating the state of the phosphorus reaction in steelmaking.

Fig . 3. Equilibrium distribution of

oxide or calcium phosphate.phosphorus between slag and iron for saturation of slag with calcium

Fig. 4. Variation of equilibrium relation KPS with tempera-

ture and iron oxide content of saturated steelmak-

ing-type slags. After Turkdogan.s~

* All the experimental data considered in this evaluation of the equilibrium relations are for slags saturated with magnesia; therefore , the independent variables used in Eq. (2) are those for the functional relation in Eq. (6) with % Fe01 as a measure of [% 0].

Review

The representation of the experimental data in the manner described did not show any effect of CaF2 on the phosphorus distribution ratio, indicating that CaF2 in moderate proportions is equivalent CaO. This subject is discussed further later in the paper in rela-tion to the ladle refining of hot metal.

For the low oxygen potentials sustained in the pres-ence of metallic iron, the sulfur dissolves in the slag as a sulfide ion. For the slag-metal reaction (1), the equilibrium relation may be represented by

0

KS- (~°-)(%FeOt) ..................(31)

where, K5 : a function of temperature and slag com-

position. Various schemes were proposed in the past to repre-sent the variation of Ks with slag composition. As suggested by the author32~ in an earlier study of the available data, the simplest form of the composition dependence of KS is that shown in Fig. 5 for wide ranges of slag compositions and temperatures of 1 530 to 1 730 °C; the relation is well represented by a single curve for the concentration of acidic oxides above 5 %. Loscher33> studied the sulfur reaction using slaps saturated with tricalcium phosphate. For melts containing about 42 % P2O5 and 3 to 6 % FeOt, re-mainder being CaO, the KS is in the range 2 to 4 which is in general accord with other data in Fig. 5. The greater scatter of the data points at low concen-trations of acidic oxides is attributed to the greater

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sensitivity of KS at very high basicities to the relative concentrations of basic oxides. For CaO-FeOt melts, the curve intercepts the ordinate at Ks 540.

Although the steelmaking methods and practices differ to varying degrees, the same fundamental prin-ciples of gas-slag-metal reactions apply to them all. In keeping with the theme of the present discourse, the deliberation here is confined to the discussion of the state of reactions during steel refining as affected by slag composition. For low phosphorus in the fur-nace charge, i.e., <0.2 % P in the hot metal, the com-

positions of slaps at the time of furnace tapping are in the following ranges : 40 to 60 % CaO, 2 to 8 % MgO, 3 to 8 % MnO, 4 to 30 % FeOt, 12 to 28 % SiO2, 1 to 3 % P2O5, 1 to 2 % A12O3, 1 to 2 % CaF2, 0.1 to 0.2 % S, and minor amounts of other oxides.

Most industrial basic slags are saturated with mag-nesia and dicalcium (magnesium) silicate and, for the low-phosphorus practice, the sum of the concentrations of CaO, FeOt, and SiO2 is in the range 87 to 90 %. In this pseudo-ternary system, the composition of the melt is that given by the univariant equilibrium for dicalcium silicate saturation at a constant tempera-ture. When the slag composition is adjusted to ac-count for entrained metallic iron and undissolved lime

particles, the total iron oxide content of the finishing slag at tap decreases with increasing concentrations of SiO2+P2O5 and dissolved CaO as indicated by the shaded areas in Fig. 6 for basic-open hearth (BOH), oxygen-top blowing (BOF), and oxygen+lime-bottom

Fig. 5.

Variation of the equilibrium constant K5 with the

concentration of acidic oxides in complex slags.

After Turkdogan.s~

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(598) Transactions ISII, Vol. 24, 1984

blowing (Q -BOP) steelmaking practices. Variations in composition within the shaded areas suggest that some basic steelmaking slags may not be saturated with dicalcium (magnesium) silicate, despite the fact that BOF and Q; BOP slags always contain entrained

particles of undissolved lime. The Q- BOP slags usu-ally contain more GaO than BOF slaps. For the range 1 550 to 1 650 °G, the tap temperature has no detectable effect on the distribution of data points in this composition diagram.

The dotted line in each diagram is for dicalcium

(magnesium) silicate saturation in the simple GaO-6%MgO-Fe0~-SiO2 melts in equilibrium with liquid iron at 1 600 °G. There is a pronounced shift in the

position of the univariant equilibrium in the composi-tion diagram for the saturated-steelmaking slags to higher concentrations of GaO and to the correspond-ing lower concentrations of Si02 as compared to the

quaternary system GaO-6%MgO-FeOt-SiO2 in equi-librium with liquid iron. This change in the melt composition to higher concentrations of lime is due to the presence of P205, MnO, Fe203, GaF2, and small amounts of other oxides which shift the saturation iso-therm to higher concentrations of GaO. Another rationale is that the slag basicity changes from low to high levels as the iron oxide content of the saturated slag increases, e.g., the V ratio is 2.0 to 2.5 for slags containing 4 to 6 % FeOt and increases to 3.4 to 3.6 for slaps containing 28 to 30 % FeOt. That is, in saturated slags the basicity is fixed by the concentra-tion of iron oxide, hence by the concentration of car-bon in the steel.

The oxygen activity in liquid steel measured with

the EMF-oxygen sensor gives the concentration of oxy-gen in solution in the steel (for low-alloy plain carbon steels). The BOF and Q; BOP data for concent-rations of oxygen in steel and total iron oxide in slag at tap, for temperatures of 1 550 and 1 650 °G and for carbon contents in the range 0.01 to 0.8 %, are distributed within the shaded area in Fig. 7(A). The dotted curve is calculated from the data on the solubility of oxygen in liquid iron and the activity of iron oxide36> along the lime-saturation isotherm for relatively simple slaps GaO(MgO)-SiO2-FeOt. The position and shape of the shaded area (where the

plant data are distributed), relative to the dotted line for the simpler saturated slaps, indicate that the activity coefficient of iron oxide in steelmaking slags decreases with increasing concentration of iron oxide and with the corresponding increase in slag basicity.

From the compositions of slag and metal samples taken at tap in BOF and Q -BOP steelmaking at tem-

peratures of 1 590 to 1 610 °G, a plot is made in Fig. 7(B) . of the product {(%MnO)/[%Mn]} (%GaO/

%Si02) against the concentration of iron oxide. The broken line drawn with the slope of six represents the slag-metal equilibrium for the manganese reaction as

given in Eq. (27). It is self evident that there is close approach to the slag-metal equilibrium with respect to the manganese reaction in oxygen-steelmaking pro-cesses, over wide ranges of carbon content of steel at tap.

In oxygen+lime bottom blowing as in Q-BOP, the slag formation begins in the melt near the tuyere zone where the oxides of iron, manganese, silicon, and phos-

phorus react rapidly with the lime powder. As the

Fig. 6. Composition of steelmaking slags containing <3 The dotted line is for Ca0-6%Mg0-Fe0-Si02

% P205 and 4 to 8 % melts saturated with

Mg0 for temperatures of 1 550 to

dicalcium silicate and magnesia.

1 650 °C;

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slag particles traverse the bath with CO bubbles, they react with carbon. Because of the secondary reaction with carbon in the bath, the concentrations of iron oxides and manganese oxide in Q- BOP slaps are lower than those in BOF (or BOP) slaps. That is, for a given concentration of carbon in the bath, the metal/slag manganese-distribution ratio in Q- BOP is higher than in BOF.

The plant data plotted in Fig. 8 show the relation between the concentrations of oxygen and carbon in steel at tap. With the exception of low-carbon melts in the Q -BOP, the oxygen contents of steel at all car-bon levels for BOF and Q-BOP steelmaking are close to the equilibrium values for a pressure of 1.5 bar CO which is an average total gas pressure in the bath. The lower oxygen contents of steel in Q-BOP at low carbon levels suggests that oxidation of CO to CO2 in the bath is well above the equilibrium values for the reaction;

CO2+C = 2CO .....................(32)

That is, when the rate of oxygen-bottom blowing ex-ceeds the rate of consumption of oxygen by carbon, manganese, and iron, the excess oxygen will oxidize the CO generated in the bath near the tuyere zone to C02, thus diluting CO and allowing the decarburiza-tion to proceed at lower levels of residual oxygen in

the steel. On the basis of the foregoing process analysis, which

was discussed elsewhere5,6~ in more detail, certain ad-vantages are anticipated in the use of argon+oxygen mixtures in the later stages of refining of steel in both the top and bottom blowing practices, particularly for steels containing less than 0.02 % C. In fact, these anticipated advantages have been realized in recent developments of combined top and bottom blowing

practices in steelmaking, e.g., LBE and K-BOP. Many previous studies of plant data have shown

that at the time of furnace tapping in all types of steel-making processes, the sulfur-distribution ratio between slag and metal is close to the equilibrium relation shown in Fig. 5.

The BOF and Q-BOP plant data for temperatures of about 1 600 °C are plotted in Fig. 9(A) showing the variation of log K 5 with the total iron oxide con-tent of the slag. For low P205 contents in slags satu-rated with both dicalcium silicate and magnesia, the concentrations of iron oxide and silica or lime in the slag are interrelated. Therefore, for doubly saturated slaps, the phosphorus-distribution ratio alone will be a single function of the concentration of iron oxide. The equilibrium relation for doubly saturated, low-

phosphorus slags is shown by the dotted curve in Fig. 9(B). The points for Q; BOP are scattered about the

Fig. 7(A). Oxygen contents of steel content of slag at tap.

is related to iron oxide

Fig. 7(B). Equilibrium relation for

(- - - -) is compared with Turkdogan.s~

manganese

plant data.

reaction

After

Review

(600) Transactions ISIJ, Vol. 24, 1964

dotted curve representing the equilibrium relation. The data for BOF, for slags containing more than about 15 % Fe01, show departures from equilibrium in favor of high phosphorus in the steel. The same

plant data are plotted in Fig. 10 showing the variation of the phosphorus-distribution ratio with the carbon

content of the steel at tap. For steels containing 0.03 to 0.05 % C at turndown, the slag/metal phosphorus distribution ratio for Q -BOP is about twice that for BOF. Furthermore, lower levels of carbon attainable by oxygen-bottom blowing brings about more exten-sive dephosphorization of the metal.

Fig. 8.

Relation between

carbon in liquid

blowing practices.

the concentrations of oxygen and

steel for oxygen-top and bottom

After Turkdogan.6~

Fig. 9. State of phosphorus reaction at time of furnace tap-

ping in BOF and Q -BOP for temperatures of about 1 600 °C is compared to the equilibrium relation

(- - - -) for saturated slags. After Turkdogan.6

Review

Transactions ISIJ, Vol. 24, 1984 (601)

As was noted in previous studies,5,6~ the analysis of the plant data substantiates the view that at the time of furnace tapping there is close approach to slag-metal equilibrium for most reactions. Therefore, the equilibrium data for slag-metal reactions in conjunc-tion with plant data will provide a reliable basis for static and dynamic process control in steelmaking, with certain adjustments in the charge control model to accommodate small differences in melt-shop prac-tices.

III. Ladle Refining

Prior to the 1970's, the ladle treatment of hot metal was confined almost entirely to desulfurization when it was considered necessary. In the case of liquid steel, the ladle refining has been practiced for a much longer time. However, prior to the 1950's, the ladle refining of steel meant deoxidation, recarburization, resulfurization for free machinability of steel and mod-erate additions of alloying elements. The advent of continuous casting and an increase in customer de-mand for high-strength, low-alloy steels with improved transverse mechanical properties made it necessary to develop increasingly more sophisticated methods of treating steel in the ladle for deoxidation, desulfuriza-tion, degassing, and shape control of inclusions. In recent years, there has been a growing demand for cleaner steels, i.e., steels containing very low concen-trations of nonmetallic inclusions, low sulfur steels (< 10 ppm S), alloy steels low in interstitial elements H, N, C, 0, and S, and for other quality steels for

special applications. Such high levels of steel refine-ment can be achieved only by a suitable method of ladle treatment.

Since the late 1970's, there has been a rapid growth of technology for ladle refining of both hot metal and steel, the so-called secondary steelmaking or ladle metallurgy. There is an overwhelming volume of

publications of research and development on ladle metallurgy. Regretably, many of the publications, both from the academia and the industry, are on re-

petitive work. A critical assessment of the accumu-lated knowledge and technical information on ladle metallurgy in a comprehensive, yet in a condensed form, would be a worthwhile undertaking. In the

present discussion of a few selected areas of ladle re-fining, with particular emphasis on slag-metal reac-tions, no attempt is made to cite in a chronological order the large volume of repetitive publications on the subject. References to almost all the publica-tions related to ladle metallurgy can be found in the

proceedings of three series of symposia, " Scaninject ", held in Lulea, Sweden, in 1977, 1980, and 1983,3'-39) and in other publications cited in this paper.

1. Rate of Refining in Stirred Melts

To enhance the rate of refining of steel with the overlaying slag, with the injected powder reactants and the flotation of reaction products out of the steel bath, various methods of stirring have been developed in the steel industry; electromagnetic, mechanical, and argon injection. References may be made to re-

Fig. 10. Phosphorus distribution bon content of steel at and Q-BOP steelmaking.

ratio

tap

is related to the car-

(4600°C) in BOF

Review

(602) Transactions ISIJ, Vol. 24, 1984

view papers40> on the practical merits of these pro-cesses. Reference may be made also to Szekely41 for

studies of better understanding of the fluidynamics of mixing in stirred melts through mathematical model-

ing and computer-aided calculations. For the pres-ent, it will suffice to state that the experimental and

plant studies have shown that the rate of slag-metal reactions or flotation of reaction products in stirred

melts can be represented by an equation of the form42~ :

Ct = C0 exp (-kt) (33)

where, C0: the initial concentration of reactant in the steel

Ct : the concentration at time t k : the apparent volumetric rate constant for

a given type and intensity of stirring. The mixing time, z, to achieve about 95 % homo-

genization in stirred melts is a good measure of the rate of ladle refining. From measurements of the rate of solute homogenization in 50 t argon stirred melts in the ladle, 50 t ASEA-SKF, 200 t RH and 65 kg water model experiments, Nakanishi et a1.43~ derived the fol-lowing empirical relation for mixing time z as a func-tion of the rate of dissipation of energy, ~, in the stirred melt.

r = (6O0±100)°° (34)

where z is in second and in W/t liquid steel. For gas stirred melts, is that due essentially to the buoy-ancy energy of the injected gas, the contributions of the kinetic energy of the gas, and the thermal expan-sion of gas to the circulating flow are small enough to be neglected.

E = 6.18 X 10-3VT In 1 +10pg5~ .........(35} W P

where, V: the gas flow rate (nm3/min) T: the melt temperature (K)

W: the mass of the melt (t) p : the density of the melt (kg/m3)

g: the gravity acceleration (m/s2) H: the depth of the gas inlet (m)

P: the gas pressure at inlet (bar). For p=7 000 kg/m3 and g=9.81 m/s2, is given by

E=

0.014VTlog 11+

H

W 1.46P(36)

Mori and Sano44~ proposed the following equation

for time of mixing for the purpose of scale-up

z=FL(W/p)2/3 1/3

(37)

where, F (dimensionless) : a proportionality factor that depends on the geometry of the melt container.

In model experiments of solute homogenization in aqueous solutions, Haida and Brimacombe45> used torpedo- and ladle-shaped vessels, and found that F increases with an increase in the diameter/height ratio, D/H, for the cylindrical ladle and length/height ratio, L/H, for the torpedo ladle. The following relation

Review

may be obtained from their results

D 0.85 L 0.85

F=150(H) =150(H) (37-a)

Haida and Brimacombe45~ also found that their re-sults on mixing time in aqueous solutions as repre-sented by Eq. (37), agree well with those of Nakanishi et a1.43~ for stirred liquid steel mentioned earlier. An-other formulation of the mixing time given below is that derived by Sano and Mori46~ from a mathematical model that describes circulating flow in a molten steel bath with inert gas injection.

z = 100(D2/H)2 ]0.337 - - ---- (38)

This derived formula is also found to agree well with the measured values of mixing time in stirred steel

baths and in water-model experiments. Kikuchi et a1.47~ measured the rate of dephosphori-

zation of low-carbon steels in stirred melts. They found that the volumetric rate constant, k', for de-

phosphorization increased with an increase in the in-tensity of stirring. Their results may be approxi-

mated by the relation:

k' (min-1) ^' 0.00 70.4 (39)

which, in combination with Eq. (34), gives k'z ~ 0.07. That is, as one would expect from the rate of mass transfer considerations, the volumetric rate constant is inversely proportional to the mixing time.

2. Hot-metal Refining

Since the late 1970's there has been an intensified R & D effort, particularly in Japan, to explore the uses of fluxes in the refining of molten pig iron to low levels of silicon, sulfur, and phosphorus. The objec-tives-with an economic and environmental bias-are two-fold :

(i) to increase the recycling of steelmaking slags (low in phosphorus) to the blast furnace for the re-covery of iron and manganese; and

(ii) to develop slagless steelmaking from hot-metal low in sulfur, phosphorus, and silicon, thus minimizing slag disposal. In his 1981 Howe Memorial Lecture, Fuwa48~ present-ed a detailed account of research and development in Japan on the ladle treatment of hot metal. This subject will not be pursued further here with the ex-ception of citing a few selected interesting examples of slag-metal reactions pertaining to the desulfurization and dephosphorization of hot metal by ladle treat-ment. In practice, the hot metal is first desiliconized to about 0.1 % Si in the tap-torpedo ladle by the addition of iron ore or manganese ore followed by nitrogen stir of the melt. After deslagging, the hot metal is desulfurized and dephosphorized by injecting with nitrogen and/or oxygen a mixture of sinter fines, burnt lime, calcium fluoride, and some calcium chlo-ride. In some practices, sodium carbonate alone is injected. At low temperatures and with fluxes having high capacities for phosphorus and sulfur, the hot

metal can be refined to the residual levels of 0.005

% P and 0.002 % S. It is well to remember that the hot metal thus refined contains hardly any silicon, therefore there is severe limitation to the amount of scrap steel that can be melted during decarburization in the converter.

At sufficiently low oxygen potentials, the phospho-rus dissolves in molten slags and fluxes as a phosphide ion (P3-). Tabuchi et a1.49) studied the following phosphide and phosphate reactions in CaO-A1203 and CaO-CaF2 melts.

--P2 + 3 (02-) = (P3~)+ 3 ... ...............(40) 2 2 4

±p24(Q2-)Q2 -E- 5 = (POr).............(41) 2 2 4

They found that the critical partial pressure of oxygen for transition from the phosphide to phosphate reac-tion at 1 550 °C is about 10-18 bar 02 for the calcium aluminate melt and about 10-16 bar 02 for the lime-saturated CaO-CaF2 melt. Their experimental data

give for the phosphate capacity, in units of wt%/ bar5/8, at 1 550 °C:

C _ (%P205) (2) / / ..................4P ` (PP2)12(Po2)54

GP=4.1 x 1022 for the aluminate melt (41 %CaO+ 59%A1203) and C =1.4 x 1021 for the lime-saturated CaO-CaF2 melt. To facilitate comparison with other slag-metal equilibria for the phosphate reaction, Cp may be converted to another formulation of the phos-

phate capacity

(%P205) Kp [%P][%0]2.5

(~ ) 1/2 P2 02 = CP {%P} [%]2 .5 ...............(43)

With the known free energy data for solution of gase-ous oxygen in liquid iron50) and of gaseous phosphorus (P2) in liquid iron,51) the CP values of Tabuchi et al.49) for 1 550 °C give KP=4.1 x 109 for the aluminate melt

(N 40 % CaO) and KP =1.4 x 108 for the fluoride melt (N20 % CaO).

In the system CaO-Fe0r-P205 saturated with CaO and Ca4P209, the melt contains about 28 % P205 and 15 % FeOt. For this saturated slag, the equilibrium constant KP in Fq. (28) gives K=2.1 x 106 for 1 550 °C and KP=0.6 x 106 for 1 600 °C. The values of KP derived from the results of Tabuchi et al.49) are two to three orders of magnitude higher than that for the lime- and phosphate-saturated melt.

In laboratory experiments, Kawai et al.52) mea-sured the concentrations of oxygen and phosphorus in iron samples equilibrated with lime/fluoride- and lime/ alumina-based slags at 1 600 °C. Their data give values of KP increasing from about 105 at 50 % (CaO +CaF2) to about 106 at 85 % (CaO+CaF2); for the aluminate melts (with no CaF2) containing 35 to 40 % CaO, the value of KP is in the range 4 x 104 to 1.3 x 105. High values of KP derived from the data of Tabuchi et a1.49) are inconsistent with other slag-metal equilibrium data. This disparity may be due

Transactions ISIJ, Vol. 24, 1984 ( 603 )

to the indirect method of estimating the oxygen poten-tial in their experimental system.

Because of current interest on fluoride-based slaps or fluxes for hot metal dephosphorization, numerous speculations have been made of the possible effect of CaF2 on the phosphate capacity of slags. An empiri-cal equation proposed by Healy29) to describe the phosphorus distribution ratio as a function of slag composition when used for fluoride slaps gives some unusual results. As shown by Kawai et a1.,52) for ex-ample, such an exercise would predict a distribution ratio (%P)/[%P] in an 80 % CaF2-based slag about four orders of magnitude greater than that in the CaF2-free slag, an impossible situation.

As discussed elsewhere,6) and briefly mentioned earlier, for lime- and magnesia-saturated slags the KPS in Eq. (29) for a given temperature is a single function of the concentration of iron oxide in the slag. On the basis of this reasoning, the data of Kawai et a1.52) and Kikuchi et a1.47) for 1 600 °C are plotted in Fig. 11 where the dotted curve represents the equi-librium relation for complex steelmaking-type slags

(containing 0 to 5 % CaF2) assessed in a previous study.6) It should be noted that Kikuchi et a1.47) gave their results in a graphical form as isoplethes of (%P)/

[%P] in a pseudo-ternary system CaO-CaF2 FeOt (total being 88 %) of MgO-saturated slags containing about 4 % Si02 and 1 to 2 % P205. Considering the diversity in slag compositions, there is a general agree-ment between the results of vastly different types of experimental work.

The alumi.nate anions copolymerize with silicate and phosphate anions to form complex alumino-silico-

phosphate anions. Therefore, in the functional rela-tion KPS we may include alumina. Assuming that 0.5 mol A1203 is equivalent to 1 mol Si02, the concen-tration of A1203 as molar silica equivalence would be 1.18 x %A1203. In the lower diagram in Fig. 11, log

(1 + %Si02+ 1.18 x %A1203) (%P205)/[%P] is plot-ted against the total iron oxide content of the slag. It follows from the foregoing analysis that CaF2 has no

particular effect on the phosphate capacity of the slag where CaF2 behaves much like CaO. The major role of CaF2 in slaps for hot metal dephosphorization is to lower the liquidus temperature and lower the viscos-ity of the slag for use at hot metal temperatures. The highest distribution ratio (%P)/[%P] attainable is with lime-saturated slags containing 15 to 25 % FeO, and low concentrations of Si02 and A1203.

As would be anticipated from the thermodynamic

properties of phosphate melts, the ratio (%P)/[%P] obtained with soda-based slaps is higher than that with lime-based slaps. For example, Marukawa et a1.53) have found that in the dephosphorization of liquid iron, essentially saturated with graphite, the ratio

(%P205)/[%P] is about 1 000 at 1 300 °C for slag basicity of Na2O/Si02=3; a similar distribution ratio is obtained at a lower basicity of 1.5 when the tem-

perature is decreased to 1 200 °C.54) It should be noted, however, that carbon increases the activity coefficient of phosphorus dissolved in iron; for hot met-al compositions the activity coefficient of phosphorus

Review

(604) Transactions ISIJ, Vol. 24, 1984

is about 5,55) which in part accounts for the attain-ment of high values of the ratio (%P)/[%P] in the dephosphorization of hot metal.

In a recent publication, Takeuchi et a1.56j presented an interesting account of transitory reactions that oc-cur in hot metal refining by injection of flux with oxy-

gen, based on their experiments with 100 t hot metal. A mixture of lime, iron ore, and fluorspar was injected with a 4: 1 02: N2 mixture at a depth of 1.5 m below the melt surface in a transfer ladle. The oxygen ac-tivity in various locations in the metal bath was mea-sured with an EMF-oxygen sensor. The oxygen ac-tivity in the bath at locations away from the gas stream and well below the slag layer were close to the value expected for the carbon-oxygen equilibrium with C0; the oxygen activity was about 100 times greater near the slag layer and 1 000 times greater near the lance tip. The variation in oxygen activity with location in the melt was in general accord with

the pattern of liquid flow. The metal was dephos-

phorized at 1 300 °C from about 0.1 to 0.01 % P with carbon remaining at about 4.4 %. From the com-

positions of metal and slag samples and measured oxy-gen activities near the slag layer, they estimated the phosphate capacity of the slag. For CaF2-fluxed slaps with basicity V N 3, they obtained values in the range 1025 to 1026 for Cp at 1 300 °C; this corresponds to Kp 1010 which is close to the value estimated by extrapolating the data for steelmaking slags from 1 600 to 1 300 °C, with due adjustment of Kpo for the ac-tivity coefficient fp N 5 in graphite saturated iron.

Compared to dephosphorization, it is much easier to desulfurize hot metal by various means, e.g., injec-tion of magnesium, calcium carbide, burnt lime, and fluoride- or soda-based slags used for dephosphori-zation. In a recent publication, Turkdogan and Martonik57~ presented some equilibrium data for the

gas-slag-metal reaction:

CaO(s)+S+C = CaS(s)+CO............(44)

It will suffice for the present to state that at 1 bar

pressure of CO and for graphite-saturated melts coex-istent with solid CaO and CaS, the equilibrium con-centration of sulfur in liquid iron is about 4 ppm at 1 600 °C and increases to about 8 ppm at 1 400 °C. With injection of burnt lime, the hot metal can, in practice, be desulfurized to about 20 ppm S. Emi and Iida,58~ however, were able to desulfurize hot met-al to 12 ppm S at about 1 400 °C by injection of pow-dered limestone into hot metal in the transfer ladle. Improved efficiency of desulfurization close to the equilibrium levels of sulfur is attributed to the higher reactivity of finely dispersed CaO that is formed upon in situ dissociation of injected limestone powder. An-other advantage is that the CO2 generated reacts exothermically with silicon in the hot metal, thus compensates for the heat losses.

3. Liquid Steel Refining

In the ladle refining of steel we are concerned

primarily with deoxidation, desulfurization and de-gassing of steel. Recently, the author59~ presented a critical evaluation of current knowledge on ladle de-oxidation, desulfurization, and inclusions in steel, both fundamental and practical aspects. The present dis-cussion of this subject, which does not include degas-sing, is an addendum to the author's previous paper.

Calcium aluminate-based slags play a key role in the ladle deoxidation and desulfurization of steel. To ensure the attainment of isotropic mechanical prop-erties in the steel, the sulfide inclusions that form during solidification of the steel should be in the form of evenly dispersed globules of sizes in the micron range. The presence of alumina inclusions in the steel is undesirable in continuous casting and in many applications of the steel; the calcium aluminate, as finely dispersed particles, is the preferred form of oxide inclusions in killed steels. These objectives can be fulfilled for killed steels containing less than a total of 25-ppm sulfur+oxygen by achieving a fine dispersion of minute particles of molten calcium aluminate in

Fig. 11, Variation of the equilibrium relation Kp5 with the

iron oxide content of calcium fluoride and calcium

aluminate slags is compared with saturated-steel-

making type slags at 1 600 °C,

Review

liquid steel during the ladle treatment.

In the ladle desulfurization of steel with aluminum and prefused calcium aluminate slag, the following re-action prevails.

(CaO)+ 2 AI+S = (CaS)+ 1(A1203) ......(45) 3 3

The equilibrium constant for this reaction determined experimentally by Schurmann et al.60~ for 1 600 °C is about three-fold of the value computed59~ from ther-mochemical data and reliable equilibrium data for various gas-slag-metal reactions. To resolve this dis-crepancy, Ozturk and Turkdogan61 have recently determined, under closely controlled experimental conditions, the equilibrium distribution of sulfur be-tween calcium aluminate melts and liquid iron con-taining aluminum. Also, the measurements were made of the solubility of CaS at 1 550 and 1 600 °C over the composition range of the liquid phase from the aluminate to the lime saturation. The results of these equilibrium measurements were found to be in close agreement with those computed previously59~ from other equilibrium data.

As is seen from the equilibrium relation in Fig. 12(A) for steel containing 0.04 % Al in solution, the sulfur-distribution ratio increases appreciably with increasing ratio CaO/ A1203. Obviously, the lime-saturated aluminate slag should be used for effective desulfurization of steel. The effect of aluminum con-centration in steel on the sulfur-distribution ratio is shown in Fig. 12(B) for the lime-saturated aluminate melts in which the solubility of CaS increases from

Transactions ISIJ, Vol. 24, 1984 (605)

1.8 % S at 1 500 °C to 2.2 % S at 1 650 °C. These equilibrium data indicate that the steel containing 0.03 to 0.05 % Al can be desulfurized to the levels of a few ppm S by interacting the argon-stirred liquid steel with the lime-saturated aluminate slag. In fact, in a well-controlled ladle refining practice,* it has been possible to desulfurize steel on industrial scale to levels below 10 ppm S.

If the ladle treatment of steel with calcium alumi-nate is intended to achieve cleaner steel, i.e., low in total oxygen as aluminate inclusions, without desul-furization, the ratio CaO/A1203 in the synthetic slag and the residual aluminum in the metal should be reduced as indicated by the equilibrium data. For ex-ample, for the steel to contain 0.02 % S and 0.02 % Al after the ladle treatment, the prefused calcium alumi-nate to be used should be saturated with sulfur, e.g., -'2 % S, and have a mass ratio CaO/A1203 of about 1: 1. This is in the middle of the liquid composition range; therefore, the slag has capacity to absorb alu-mina inclusions from the steel when mixed with the metal with argon stirring. For the particular case considered as an example, the equilibrium concentra-tion of oxygen in solution in the steel would be about 2 to 3 ppm.

The aluminum-killed steel may be desulfurized also by injecting a mixture of burnt lime and fluorspar. The extent of desulfurization that can be achieved depends on the efficiency of fluxing of the deoxidation

product alumina with the injected lime. Kor and Richardson62> determined the sulfide capacity, GS, of CaO-CaF2 A1203 melts at 1 500 °C. Their data are

Fig

(a) (b)

. 12. Equilibrium sulfur distribution between liquid calcium aluminate and low alloy steel.

No slag carry-over during furnace tapp

Basic refractory lining for the ladle

ing (c) (d)

High argon flow using porous plug

Ladle cover to avoid air oxidation of th

After Ozturk and Turkdogan.61

e melt

Review

( 606) Transactions ISIJ, Vol. 24, 1984

presented in a different form in Fig. 13 as a plot of log KS against the concentration of CaF2 for indicated mass ratios of %A1203/%CaO. It is seen that CaF2 has only a small effect on the sulfide capacity of the aluminate slag. However, CaF2 fluxes both CaO and CaS; for the invariant equilibrium at 1 500 °C, the liquid phase contains about 60 % CaF2, 20 % CaS, and 20 % CaO. Therefore, an addition of 5 to 10 % CaF2 to the injected burnt lime would curtail the coat-ing of lime particles with a layer of CaS, hence would improve the efficiency of lime usage in ladle desul-furization. In the silicon-deoxidized steel, the desulfurization

is governed by the reaction :

(CaO)+ Si+S = (CaS)+ 1(Si02) 2 2 ............................(46) .

The sulfide capacity, and the deoxidation capability with silicon, of calcium silicate melts is low even at lime saturation; therefore, desulfurization by reaction (46) is possible only when the activity of silica is low-ered significantly by using a suitable flux such as pre-fused calcium aluminate. The relation shown in Fig. 14 for sulfur and silicon concentrations is for steels equilibrated with calcium aluminate+5%Si02 melts saturated with CaO and CaS. In this slag, the ac-tivity of silica is in the range 10_5 to 10_4 and the activity of alumina about 5 X 1013, both with respect to the pure solid oxides, for which the following are the equilibrium concentrations of silicon and alumi-num in the steel at 1 600 °C: 0.01 % Al for 0.07 % Si and 0.03 % Al for 0.3 % Si. However, at these low oxide activities in the slag considered, the kinetics of Si-Al redox reaction would be sluggish, consequently at low solute concentrations, e.g., <0.05 % Al and <0.5 % Si, there may be little or no countercurrent transfer of silicon and aluminum between slag and metal by the redox reaction in a relatively short time of ladle refining. However, if the aluminate slag con-tains 8 to 10 % Si02, there could be appreciable loss of aluminum from metal to slag because of oxidation

of aluminum in the steel by silica in the slag.

Iv, Role of Vapor Species

Because of the high temperatures involved, vapori-zation plays a significant role in pyrometallurgical re-actions. In a reactor such as a blast furnace, the formation and condensation of various vapor species is responsible for the recycle of some of the elements between the high and low temperature regions of the furnace. Familiar iron oxide fumes emitted in steel-making operations is a consequence of enhanced va-porization of iron under certain conditions. In the nonferrous industry, the slag fuming is used in the recovery of, for example, tin and zinc from slags. Va-

porization in high-temperature manufacturing pro-cesses (i) can be an asset for the operation of the process, (ii) may impede the smooth operation of the process, or (iii) may cause pollution of the environ-ment. Many aspects of vaporization studied in the

past few decades have been of much value in the de-velopment of special processing techniques, a better understanding of the workings of the high-temperature

processes and in curtailing vaporization-related pollu-tion problems.

1. Vaporization in the Blast Furnace

The countercurrent flow in the blast furnace of

solids from low- to high-temperature zones and of

gases from high- to low-temperature zones brings about a cyclic process of vaporization and condensa-tion which has a decisive influence on the overall

operation of the furnace. It is now well established that the silicon and sulfur in the coke are transferred

to the slag and iron in the blast furnace via the vapor species SiO, SiS, CS, and other minor sulfur-bearing

species. The earlier and recent studies of quenched

Fig. 13. Sulfide capacity KS of CaO-CaF2-A120

1 500 °C derived from data of Kor and

son.62~

3 slaps at

Richard-

Fig. 14. Equilibrium sulfur and silicon solubility in liquid steel coexistent with calcium aluminate+5 % Si02 slag saturated with CaO and CaS. After Ozturk and Turkdogan.si~

Review

Transactions ISII, Vol. 24, 1984 (607)

blast furnaces have revealed that while the sulfur con-tent of the slag increases with the descent of the burden in the bosh, the sulfur and silicon contents of the metal droplets reach maxima and then decrease as they pass through the slag layer.19'63,64) This phenomenon of vapor-phase mass transfer has been substantiated by the experimental work of Tsuchiya et a1.65) and Tukr-dogan et a1.14) under conditions partly simulating the bosh region of the blast furnace.

In the presence of carbon and depending on tem-perature and activity of silica in the slag or coke, the SiO vapor is generated by one of the following reac-tions.

Si02 (in coke ash or slag)+C --~ SiO+CO......(47)

SiC+CO -> SiO+2C .................................(48)

At 1 bar pressure of CO, as1o2=1 and above about 1 500 °C, Si02 is converted to SiC and SiO is gen-erated by reaction (48); for a5102=0.1, reaction (48) applies only at temperatures above about 1 610 °C. In the experiments cited,14) the SiO generated by pas-sage of CO through a coke bed at temperatures of 1 600 to 1900 °C, were found to be close to the equi-librium values for reaction (48). Volatile SiS is gen-erated by reactions of the type with the coke ash :

CaS (in coke ash)+SiO --> SiS+CaO ............(49)

FeS (in coke ash)+SiO+C -p SiS+CO+Fe

........................(50)

The reaction of these vapor species with metal and slag may be represented by the equations:

SiO+C = Si+CO Metal SiSSi

+S ..............................(51)

=

SiO+0.5 (02-)+0.5C = ( )Si-O-)+0.5C0 Slag SiS+ 1.5 (02-)+0.5C = ( )Si-O-)+(S2-)

+0.500 ........................(52)

Experiments have been made to simulate the events in the upper part of the blast-furnace hearth where metal droplets pass through the slag layer.14) Within a short residence time (1 to 2 s) of metal droplets in the slag, appreciable changes occurred in the composi-tion of the metal droplets : manganese oxide in the slag oxidized silicon in metal droplets from 1.0 % Si to about 0.2 % Si and the metal was desulfurized from 0.1 % S to about 0.02 % S.

Coke samples taken from the tuyere zone of the blast furnace usually contain 12 to 16 % ash which has a basicity (CaO+MgO)/Si02 of about 0.6; for a typical blast furnace slag, the basicity is about 1.5. Therefore, the partial pressures of SiO and sulfur-bearing species in the gas near the coke surface are expected to differ from those interacting with the slag. For the purpose of comparison, calculations are made for assumed local equilibrium at 1 500 °C for the sys-tems gas-slag and gas-coke ash; the results are given in Table 1. The amount of ash in the tuyere coke and its composition suggests that, although most of the silica and sulfur in the coke ash are removed from

the furnace by the slag and liquid iron, there is sig-nificant recycle of silicon and sulfur in the bosh region by the vaporization and condensation of these species on the coke particles.

The alkalies constitute another dominant vapor spe-cies in the bosh and stack regions of the blast furnace. Many studies have been made of the alkali recycle in the blast furnace via the process of vaporization and condensation; the subject is well covered in the ref-erences cited.66,67) General concensus evolved from elementary thermodynamic calculations is that the alkali silicates in the ore and coke ash decompose at elevated temperatures in the lower part of the bosh and the combustion zone. In part the alkali vapors carried away with the ascending gas react with the slag and can therefore be removed from the furnace; part are converted to alkali cyanides and carbonates and deposited on the burden and the refractory lining of the furnace stack. Some new thoughts are pre-sented in the following discussion on the sequence of reactions and vapor species that may govern the alkali recycle in the blast furnace.

As an example, let us calculate the equilibrium va-

por pressure of potassium for the reduction reaction :

(K20)+C = 2K (g)+CO (g) ............(53)

The computed equilibrium values in Table 2 are for a blast furnace slag containing 0.5 % K20 and having a basicity of B'= 1.5 : K is the equilibrium constant for reaction (53) from the thermochemical data and aK2o is the activity of K20 in the slag, relative to the hypothetical solid K20, from the experimental data of Steiler.6S)

If the potassium input with the blast furnace burden is 4 kg. Kf t-HM and 80 % of it is removed by the slag, about 0.8 kg. Kit-HM will be in the vapor phase

Table 1. Calculated equilibrium vapor pressures in

bar for the systems gas-slag and gas-coke

ash at 1 500 °C and 1 bar pressure of CO.

Table 2. Equilibrium reaction (53)

and B' -1.5.

vapor

with

pressure of potassium for slag having 0.5 % K20

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(608) Transactions ISIJ, Vol. 24, 1984

carried away by the furnace gas. For a typical total

gas volume of 2 300 nm3/t-HM, the partial pressure of the total potassium vapor species would be 7.2 X 10.4 bar (for a total gas pressure of 3.6 bar in the bosh region). Comparison of this total potassium pressure in the gas with those in Table 2 for the gasslag equi-librium indicates that the slag particles will pick up

potassium from the gas only at temperatures below 1 200 °C. The ash in the coke samples taken from the tuyere zone contains 2 to 4 % (Na20+K20) which is several fold of that in the original coke ash. Because of the low basicity of coke ash, the amount of alkali therein is expected to be 5 to 10 times greater than that found in the slag. In the upper part of the bosh, the slag and coke will pick up alkalies from the

gas. As the temperature of the slag and coke increases during descent in the bosh, they will emit alkali vapors to the gas phase. However, the partial pressure of potassium vapor in the gas would probably be below that given in Table 2 for the gas-slag equilibrium. Nevertheless, the numerical examples cited suggest that much of the alkalies accumulated in the bosh and hearth zone of the blast furnace, via recycling, are in the gas phase.

According to a mass spectrometric study of vapori-zation of potassium cyanide by Simmons et al.,69) the equilibrium pressures of vapor species are in the order p2 (KCN) > PKCN > P3 (KCN) . For atmospheric pressure of nitrogen and at graphite saturation, the following are the equilibrium ratios of the vapor species : at the melting point of KCN (635 °C), (KCN)2: KCN: K =2: 1: 0.76 and at the boiling point (1 132 °C),

(KCN)2: KCN: K=4.6: 1: 0.19.* It is all too clear that (KCN)2 is the dominant alkali-bearing species in the blast furnace. The sodium in the burden under-

goes cyclic reactions similarly to potassium; however, the concentration of sodium-bearing species are about one-tenth of the potassium-bearing species.

In addition to the well-known problems associated with alkali accumulation in the blast furnace stack, the alkali cyanides are also responsible for the genera-tion of ammonia and hydrogen cyanide in the upper

part of the stack by the reactions :

2KCN+3H20 = K2C03+2NH3+C .....

NH3+C0 = HCN+H20 ....................

.(54)

(55)

These and other organic volatile matters thus gen-erated in the stack end up in the wash water of the blast furnace off-gas. Recent experimental work of Turkdogan and Josephic,70) under conditions partly simulating the blast furnace stack, substantiated the validity of this mechanism of generation of ammonia and hydrogen cyanide in the stack.

2. Iron Oxide Fumes

Evolution of iron oxide fumes is a common occur-

rence in all steelmaking operations, e.g., in melting, in refining with oxygen blowing, and during furnace tap-

ping. The mechanism of this interesting phenom-enon was resolved in an earlier study by Turkdogan et al.,71) based on the concept of diffusion-limited en-hanced vaporization of metals in an oxidizing envir-onment. In this counterflux-transport process, at some short distance from the surface of the metal, the oxygen and metal vapor react to form a metal oxide mist (fume). The formation of an oxide in the gas

phase provides a sink for the metal vapor and oxygen resulting in the counterflux of these two gaseous spe-cies. An increase in the oxygen partial pressure, and or an increase in the gas-film mass-transfer coefficient for oxygen, decreases the thickness of the sub-bound-ary layer through which the metal vapor is diffusing, hence increases the rate of fuming. Also, the higher the vapor pressure, or higher the temperature, the faster would be the rate of fuming. For given tem-

perature and conditions of gas flow over the metal sur-face, the rate of vaporization is represented in a gen-eral form by the relation :

Rate of fuming = P(1-0) exp (-EJRT)

_ 0(1- D)p~,, }

where,

(56)

When

metal diffusi

becom -f

the fur fumin zation expect p tally. During furnace tapping, the metal stream is ex-

posed to air, consequently the iron oxide fumes are emitted by the mechanism described. The fume emis-sion is particularly intensive in the blast furnace metal runner, because of the high vapor pressure of man-

ganese at the concentrations present in the hot metal. Furthermore, the metal droplets ejected by gas evolu-tion from the hot metal in the runner increase the

gas-metal interfacial area, consequently intensify the fume emission. It follows from the basic concept of the mechanism of fuming that the extent of fume emis-sion can be suppressed by either eliminating gaseous oxygen from the exposed surface of the metal or by covering the metal surface with a layer of slag. De-

pletion of oxygen near the metal surface can be achieved by shrouding the metal runner with a gas

F: is the energy of vaporization b, ~ : constants for given conditions of

steady-state vaporization

(1-0) : the fraction of metal surface exposed to the oxidizing gas stream.

the flux of oxygen toward the surface of the is greater than the equivalent countercurrent

ve flux of the metal vapor, the metal surface es oxidized, i.e., (1-D) 0, and consequently

ing ceases. Just before this cutoff, the rate of g will be close to the maximum rate of vapori-that is attainable is vacuo. These theoretical ations have been well substantiated ex erimen-

* It should be noted that in the free energy tables in p . 12 of Ref. 5), the value given for the vaporization of liquid KCN is incorrect; it should have been, according to the data of Simmons et al.ss~

2KCN (1) _ (KCN)2 (g) 4G° = 52180-36.64Y(cal)

Review

flame. Extensive plant trials done recently in U.S. Steel have substantiated the viability of this method

of fume suppression during furnace tapping.

3. White Fumes from Blast Furnace Slag Runner

White fumes emitted from the blast furnace slag runner is another example of vaporization-induced oc-currence. Samples of white fumes collected over the slag runner have been characterized recently by Simon and Kelly,72~ in U.S. Steel Research Laboratory, using the scanning electron microscope. The X-ray scan spectra revealed that the potassium and sulfur were the most dominant elements in the fume samples. Such a finding is consistent with the expected reaction of slag with air during tapping.

The blast furnace slag contains 1.5 to 2.0 % S, and as discussed earlier, the potassium vapor pressure of the slag is relatively high. Therefore, upon exposer to air, the potassium and sulfur will be evolved from the slag as potassium sulfate by the reaction:

2 (K+)+(S2-)+202 = K2SO4 (white fume) ...........................(57)

In addition, there is SO2 evolution, readily evidenced by the characteristic odor in the casthouse,

(S2-)+ 3 02 ` (02-)+S02 ............(58) 2

In the oxidizing atmosphere over the slag, the sulfur dioxide is partly oxidized to sulfur trioxide. In the cooler ambient atmosphere away from the slag sur-face, 503 and K2SO4 will react with moisture in the air and form a sulfate complex xK2SO4.yH2SO4 which appears as white fume. The X-ray scan spectra indi-cated an atom concentration ratio K/S of about one in the collected fume samples, which apparently has the composition of K2SO4. H2SO4. The higher the slag temperature, the greater the extent of fuming caused by the above reactions.

The emission of white fumes can be suppressed by spreading a thin layer of sand over the slag surface in the runner. This simple remedy works for two rea-sons : added sand, or other inert powder, chills the salg surface and also reduces the surface area of slag ex-posed to air.

Yamamoto and Harashima73~ studied changes in chemical composition of slag during the refining of hot metal with sodium carbonate and oxygen-top blowing. They deduced from the composition of slag and fume collected that there was preferential vapori-zation of sulfur and sodium from the slag. This ob-servation can be explained also in terms of reaction (57) written for Na+ in the slag and Na2SO4 in the fume.

4. Vaporization in Coal Gas cation

Certain vapor species that are generated in some

coal gasification processes may cause side reactions which could affect the industrial application of the

gasification process. Because such processes are still in the early stages of development, technical informa-

Transactions ISIJ, Vol. 24, 1984 ( 609)

tion available is meagre, therefore we can at present only speculate on the possibility of vaporization-in-duced side reactions and consequences thereof.

In the coal gasification processes that are now being developed, e.g., CGS, COIN, KR, the pulverized or

granular coal is partially combusted with oxygen and steam in a bath of liquid iron to generate a reducing gas for use in the ore reduction and smelting processes. In his doctorate dissertation, Fukagawa74~ presented a comprehensive review of much of the work done to date on coal gasification using liquid iron as the re-action medium. Technical developments made to date on direct usage of coal in the steel industry are

presented in a series of papers published recently in Ironmaking and Steelmaking.75~

Bench scale experiments and pilot plant tests have shown that about 15 to 30 % of the sulfur in coal is absorbed by the molten iron and slag in the gasifier and the remainder transferred to the gas and iron dust

(fume) emitted. The gas--metal interaction in the combustion zone is conducive to enhanced vaporiza-tion of iron by the mechanism already described. Away from the combustion zone, the iron oxide fume is reduced to metallic iron dust which interacts with sulfur-bearing vapor species in the reducing gas. As the temperature of the dust-laden gas decreases, more sulfur is absorbed by the iron dust. Upon cooling and dedusting, the reducing gas generated contains less than 300 vol-ppm H2S. The iron oxide fume emitted during gasification of coal is an efficient gas desulfurizer contained in the system only when the reducing gas is cooled and dedusted. The dust re-moved from the cooled reducing gas contains up to about 8 % S, a material not suitable for recycle.

The sulfur and dust-related problems with these coal gasifiers have to be surmounted for direct usage of hot reducing gas in the shaft or fluidized bed reduc-tion of iron ore. Noting that the hot reducing gas enters the ore reducer at about 900 °C, the gas has to be dedusted at a relatively high temperature, for which a new technology is needed. In a reducing

gas containing 25 vol% H2, the entrained iron dust will not desulfurize the gas below 800 vol-ppm H2S at 900 °C. Such a high sulfur-bearing gas is not suitable for the reduction of iron ore. Hot dedusting and de-sulfurization may be achieved by passing the gas, for example, through a bed of calcined dolomite at about 950 °C. In such a process, the sulfided doloma has to be removed periodically from the reactor and be treated for disposal. Of course, other methods of hot dedusting and desulfurization can be visualized; the subject will not be pursued further here.

Another problem to be resolved is that involving alkalies in the coal. Since coal contains some nitro-

gen, most of the alkalies will be contained in the gas generated as alkali cyanide vapors, which will even-tually condense in the ore reducer. Obviously, mea-

sures have to be taken to prevent the accumulation of alkali carbonates and cyanides in the upper part of

the shaft reducer. We see from the examples cited, there are many

ramifications to the role of vapor species in pyrometal-

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(610) Transactions ISIJ, Vol. 24, 1984

lurgical or other high temperature processes. The

vapor species generated may cause pollution prob-

lems, may impede the operation of the process or may

assist the process reactions.

V. Concluding Remarks

Scope of the physical chemistry of ironmaking and steelmaking is, of course, much greater than what has been presented in this discourse. Therefore, the con-cluding remarks are confined to the subjects discussed in this paper. The state of reactions in pyrometal-lurgical processes can be evaluated with confidence from accumulated knowledge of the physicochemical and thermodynamic properties of materials and of the process engineering concepts pertinent to high temperature technology. The analysis of plant data indicates departures from equilibrium for various

gas-slag-metal reactions in the bosh and hearth zones of the blast furnace. The least departure from equilibrium is for the distribution of silicon between metal and slag. Departure from equilibrium for the manganese-silicon reaction is greater for slags of low basicity. Despite departures from equilibrium, the silicon and sulfur contents of the hot metal change in a systematic manner for a given blast furnace, i.e., high silicon/low sulfur and low silicon/high sulfur.

In oxygen steelmaking processes, most reactions ap-

proach slag-metal equilibrium at the time of furnace tapping. However, some differences are seen in the behavior of phosphorus reaction in the oxygen-top and -bottom blowing processes. The salg/metal phospho-rus distribution is closer to equilibrium in Q; BOP than in BOF steelmaking, particularly for low carbon con-tents at tap.

A greater use is being made of the gas-slag-metal equilibrium data in recent developments of ladle re-fining of hot metal and steel. At temperatures of about 1 300 °C and with fluxed-basic slags of low melting temperatures, it is now possible to dephospho-rize hot metal to the levels of 0.005 % P at carbon contents in the range 4.0 to 4.5 %. The aluminum-killed steel can be desulfurized to levels below 10 ppm S by interacting liquid steel with a lime-saturated aluminate slag using argon stirring for slag-metal mixing. Vapor species are now known to play various roles in pyrometallurgical processes. The alkali cyanide vapors are responsible for alkali recycle and accumula-tion in the blast furnace and also for the generation of ammonia and hydrogen cyanide in the upper part of the stack. The silicon and sulfur-bearing vapor spe-cies are responsible for the transfer of silicon and sul-fur from coke to metal and slag in the bosh region of the blast furnace. Fumes emitted upon exposure of liquid steel, hot metal or blast furnace slag to air are consequences of some vaporization phenomenon.

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Review

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* The material in this paper is intended for general information only. Any use of this material in relation to any specific application

should be based on independent examination and verification of its unrestricted availability for such use, and a determination of suit-

ability for the application by professionally qualified personnel. No license under any United States Steel Corporation patents or other

proprietary interest is implied by the publication of this paper. Those making use of or relying upon the material assume all risks and liability arising from such use or reliance.

Review